TítuloNOx Reduction in Diesel Hydrogen Engines Using Different Strategies of Ammonia Injection

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(1)energies Article. NOx Reduction in Diesel-Hydrogen Engines Using Different Strategies of Ammonia Injection M. I. Lamas 1, * 1 2. *. and C. G. Rodriguez 2. Nautical Sciences and Marine Engineering Department, Universidade da Coruña, 15403 Ferrol, Spain Norplan Engineering S.L., 15570 Naron, Spain; [email protected] Correspondence: [email protected]; Tel.: +34-881013896. Received: 19 February 2019; Accepted: 28 March 2019; Published: 1 April 2019. . Abstract: In order to reduce NOx emissions in internal combustion engines, the present work analyzes a measurement which consists of injecting ammonia directly into the combustion chamber. A commercial compression ignition engine fueled with a hydrogen-diesel blend was studied numerically. It was verified that the flow rate shape in which the ammonia was injected, particularly rectangular, triangular, or parabolic, as well as the injection duration had an important influence on NOx reduction. A 11.4% improvement in NOx reduction, corresponding to an overall reduction of 78.2% in NOx , was found for parabolic injection shape and 1º injection duration. The effect on carbon dioxide, carbon monoxide, and hydrocarbon emissions, as well as brake-specific consumption, was negligible. Keywords: CFD; NOx ; hydrogen; ammonia; engine. 1. Introduction Nowadays, it is very important to reduce emissions in internal combustion engines as well as the dependency on fossil fuels. The current concerns about the limitation of fossil fuels and pollution are strong incentives for finding alternative sources. In this regard, both hydrogen and ammonia constitute promising fuels because these may be produced from alternative sources [1]. Since ammonia and hydrogen are carbon-free, they do not produce carbon dioxide (CO2 ), carbon monoxide (CO), unburned hydrocarbons (HC), nor particulates. Ammonia and hydrogen are also sulphur-free, and consequently, the emissions of sulphur oxides (SOx ) are also negligible, since these depend on the sulphur content in the fuel [2–5]. In internal combustion engines, ammonia-hydrogen blends complement their properties since ammonia is characterized by a narrow flammability range, high ignition energy, low combustion velocity, and high auto-ignition temperature. On the other hand, hydrogen presents a wide flammability range, low ignition energy, high combustion velocity, and lower auto-ignition temperature [6,7]. Besides, hydrogen can be obtained directly from ammonia [8–11]. Some investigations [12–15] proposed hydrogen as a promoter to speed up diesel-ammonia combustion. Technologically, the most employed option is to inject ammonia into the intake air. Boretti [16] analyzed several injection procedures in a dual fuel ammonia-hydrogen engine and concluded that the most adequate option is to inject ammonia into the intake air. Sahin et al. [17] analyzed ammonia injection into the intake air of a small engine and obtained advantages regarding efficiency and CO2 emissions, but at expenses of increments in NOx , HC, and CO. Reiter et al. [18] also employed injection of ammonia into the intake air of an engine and obtained a noticeable reduction in soot emissions but at the expense of increments in NOx , HC, CO, and consumption. The authors indicated that the increment in NOx emissions is mainly produced by ammonia oxidation. Nevertheless, they obtained that when less than 40% of the fuel energy is provided by ammonia, NOx emissions are lower than for a diesel engine because the substitution by ammonia lowers the combustion temperature and thus reduces Energies 2019, 12, 1255; doi:10.3390/en12071255. www.mdpi.com/journal/energies.

(2) Energies 2019, 12, 1255. 2 of 13. the thermal NOx formation mechanism. Another ammonia fumigation study was developed by Gill et al. [19], who injected ammonia and hydrogen into the intake air of a diesel engine and obtained reductions in consumption and CO2 emissions as well as increments in HC and NOx emissions. Hydrogen usually increases NOx emissions, especially at high loads, due to the high temperatures reached in the combustion chamber [20–22]. This constitutes an important problem, since the transport sector, especially internal combustion engines, is considered the major source of NOx formation worldwide [23,24]. The production of NOx from oxygen and nitrogen in the air is a problem associated to the combustion of fuels, and its reduction has been investigated for a long time. As a possible solution to reduce NOx , hydrogen allows internal combustion engines to operate under ultra-lean combustion and thus under low-temperature combustion and lower NOx emissions, but at the expense of reducing efficiency noticeably [25–27]. Other procedures to reduce NOx emissions in internal combustion engines can be briefly summarized in primary and secondary measures. The latter consist of removing NOx from the exhaust gases by downstream cleaning techniques, while the former consist of reducing NOx produced during combustion by reducing the concentrations of oxygen and nitrogen as well as controlling the temperatures reached in the combustion process. The high price of after-treatment systems has promoted primary measures. In hydrogen engines, several techniques have been documented in the literature, such as modification of the injection parameters, reduction of the compression ratio, exhaust gas recirculation (EGR), water injection, etc. As another measure to reduce NOx emissions in hydrogen-ammonia engines, several authors proposed to inject ammonia directly into the cylinder during the expansion stroke. As this is a novelty and recent technology, the literature about direct ammonia injection is still scarce. One can refer to Miyamoto et al. [28], who analyzed a diesel engine and showed that NOx emissions can be reduced by up to 60% by optimizing the injection timings. Larbi and Bessrorur [29] also found a substantial reduction in NOx emissions by injecting ammonia directly into the combustion chamber. Lamas et al. [30] verified that ammonia injection is much more efficient than water injection to reduce NOx . Against this background, it can be concluded that an extended amount of literature has been developed on the applications of ammonia and hydrogen as fuels. Nevertheless, the research about direct ammonia injection into the combustion chamber as a NOx reduction measure is still scarce. The present paper proposes a CFD (Computational Fluid Dynamics) model to analyze NOx reduction in an internal combustion engine fueled with a hydrogen-diesel blend. Particularly, this work analyzes the commercial compression ignition engine MAN D2840LE V10. The process of NOx reduction from ammonia injection directly into the combustion chamber was analyzed. Since ammonia is toxic, it is important to avoid un-reacted ammonia to slip into the atmosphere. For this reason, several injection shapes and durations were analyzed in order to optimize NOx reduction and thus reduce ammonia slip. 2. Materials and Methods The MAN D2840LE V10 analyzed, Figure 1, is a four-stroke, turbocharged, and direct injection engine. Table 1 summarizes the main technical specifications..

(3) Energies 2019, 12, 1255. 3 of 13. Energies 2019, 12, x FOR PEER REVIEW. 3 of 14. Figure analyzed in the present work. Figure1.1.Engine Engine analyzed in the present work. Table specifications at 100% load.load. Table1.1.Technical Technical specifications at 100%. Parameter. Parameter. Power (kW) Power (kW) Speed (rpm) Speed (rpm) Number of cylinders Number of cylinders Cylinder displacement volume (cm3 ) Bore (mm) Cylinder displacement volume (cm33) Stroke (mm) Bore (mm) Stroke (mm). Value Value 320. 320 1500 10 10 18,270 128 18270 142 128. 1500. 142. The photograph in Figure 2 illustrates the combustion chamber at the piston head. The The photograph in Figure 2 illustrates the combustion chamber at theinpiston head. The computational grid is shown in Figure 3. A deforming mesh was employed order to implement the computational is shown in FigureFigure 3. A deforming mesh was employed in ordergrid to implement the dead movement to thegrid piston and valves. 3a represents the tri-dimensional at the bottom movement to the3b piston and valves. 3a represents the elements tri-dimensional grid at the bottom dead center and Figure represents the Figure AA cross section. All are hexahedral, and the mesh size center and Figure 3b represents the AA cross section. All elements are hexahedral, and the mesh sizemeshes was refined around the valves in order to adapt properly to their opening and closing. Several was refined around the valves in order to adapt properly to their opening and closing. Several meshes with different elements and sizes were tested to verify the adequacy of the mesh. with different elements and sizes were tested to verify the adequacy of the mesh.. Figure 2. Piston head of the engine studied in the present work..

(4) Energies 2019, 12, x FOR PEER REVIEW. 4 of 14. Energies 2019, 12, 1255. 4 of 13. Figure 2. Piston head of the engine studied in the present work.. (a). (b). Figure 3. Computational mesh at bottom dead center. (a) Tri-dimensional view; (b) AA section. Figure 3. Computational mesh at bottom dead center. (a) Tri-dimensional view; (b) AA section.. The open software OpenFOAM was employed for the simulations. A new in-house solver was employed for the simulations. new in-house solver based basedThe onopen C++ software languageOpenFOAM was programmed. The solver is based on theARANS (Reynolds-averaged on C++ language was programmed. solver is based and on energy, the RANS (Reynolds-averaged Navier–Stokes) equations of conservationThe of mass, momentum, Equations (1)–(3). Navier–Stokes) equations of conservation of mass, momentum, and energy, Equations (1–3). ∂ρ ∂ (1) ∂t  + ∂x  i (ρui ) = 0  (  ui )  0 (1) t xi ∂p ∂τij ∂ ∂ ∂ (ρui ) + (ρui u j ) = − + + (−ρui0 u0j ) (2) ∂t ∂x j ∂xi ∂x j ∂x j  ij   p  ' ' (  ui )  (  ui u j )     (   ui u j ) (2) ∂ µ ∂ ∂ t ∂H x t  x  x  x j (ρui H ) = i (ρH ) +j +j Srad (3) ∂t ∂xi ∂xi σt ∂xi.    Htensor, In the equations above, ρ isthe density, τ(ijisu the viscous stress σ is the turbulent Prandtl t ( H )  H)  (3)    S rad t i t xH xi   t computed xi  number, µt is the turbulent viscosity, and is the total enthalpy, by the sum of the enthalpy i of all species. The enthalpy of each species was computed by In the equations above, ρ is the density, τij is the viscous stress tensor, σt is the turbulent Prandtl Z T number, μt is the turbulent viscosity, and H is the total enthalpy, computed by the sum of the enthalpy hk = c p,k dT + h0k ( Tre f ,k ) (4) of all species. The enthalpy of each species was Tre f computed by. . T. hk  hk (Tref ,enthalpy where cp,k is the specific heat of species k, and c h0kp ,is the  formation at the reference temperature (4) k dT k) Tref Tref,k . Thecp,kterm Srad specific in Equation a sourcek,to and include heat transfer. The P1at radiation model where is the heat(3)ofis species is the formation enthalpy the reference hk0 radiation was employed, according to which the radiative source term for the enthalpy equation is: temperature Tref,k . 0. The term Srad in Equation (3) is a source to include 4radiation heat transfer. The P1 radiation S = aG − 4aσT (5) model was employed, according to which rad the radiative source term for the enthalpy equation is: where a is the absorption coefficient, 0.2 [31], G isthe radiation, obtained by the following Sradand  aG 4aincident T4 (5) transport equation: where a is the absorption coefficient, 1 0.2 [31], and G is 4the incident radiation, obtained by the ∇ · ∇ G − aG + 4aσT = 0 (6) following transport equation: 3a   A common procedure to model stresses,  the G   aG  4a T−4 ρu  i0u j , is the Boussinesq hypothesis(6)to  Reynolds 3a   relate the Reynolds stresses to the mean velocity gradients, Equation (7). 1. 0. 0. !   ui' u 'j , is the Boussinesq hypothesis to A common procedure to model the Reynolds stresses, ∂u j ∂ui 2 0 0 −mean ρui u j velocity = µt + − ρkδij (7). (7) relate the Reynolds stresses to the gradients, Equation ∂x j ∂xi 3.

(5) Energies 2019, 12, 1255. 5 of 13. where δij is the Kronecker delta (δi = 1 if i = j and δi =0 if i 6= j), which is included to make the formula applicable to the normal Reynolds stresses for which i = j, Lamas and Rodriguez [32]. The turbulent viscosity was modeled by the k-ε turbulence model, according to which µt = ρCµ k2 /ε. The k-ε turbulence model implements two additional transport equations, one for the turbulence kinetic energy, k, Equation (8) and a further one for its dissipation rate, ε, Equation (9). ! ∂ ∂ ∂ µt ∂k (ρk) + (ρkui ) = + 2µt Sij · Sij − ρε (8) ∂t ∂xi ∂x j σk ∂x j ! ∂ ∂ ∂ µt ∂ε ε ε2 (ρε) + (ρεui ) = + C1ε 2µt Sij · Sij − C2ε ρ (9) ∂t ∂xi ∂x j σε ∂x j k k In the equations above Sij is the rate of deformation tensor, C1ε and C2ε are constants and the terms σk and σε represent the turbulent Prandtl numbers for k and ε, respectively. The default values were assumed for the model constants, Cµ = 0.09, σk = 1.00, σε = 1.30, C1ε = 1.44, and C2ε = 1.92. Regarding the simulation of fuel droplet breakup, the Kelvin-Helmholtz and Rayleigh-Taylor breakup models [33] were employed, and the heat-up and evaporation of the droplets was modeled by the Dukowicz model [34]. In order to solve the chemical kinetics, several additional equations must be added to the model. Given a set of N species and m reactions, Equation (10), the local mass fraction of each species, fk , can be expressed by Equation (11). N. kj N. ∑ vkj Mk ↔ ∑ vkj Mk j = 1, 2,. k =1. 0. 00. ..., m. (10). k =1. ∂ ∂ ∂ (ρui f k ) = (ρ f k ) + ∂t ∂xi ∂xi. . µt ∂ f k Sct ∂xi. . + Sk. (11). 0. In the equations above, vkj are the stoichiometric coefficients of the reactant species Mk in the reaction 00. j, vkj is the stoichiometric coefficients of the product species Mk in the reaction j, Sct is the turbulent Smidth number, and Sk is the net rate of production of the species Mk by chemical reaction, given by the molecular weight multiplied by the production rate of the species, Equation (12), Sk = MWk. d [ Mk ] dt. (12). where MWk is the molecular weight of the species Mk and [Mk ] its concentration. The net progress rate is given by the production of the species Mk minus the destruction of the species Mk along the m reactions ( " #) m N N 0 00 d [ Mk ] vkj vkj 00 0 = ∑ (vkj − vkj ) k f j ∏ [ Mk ] − k bj ∏ [ Mk ] (13) dt j =1 k =1 k =1 where kfj and kfb are the forward and backward reaction rate constants for each reaction j. A reaction mechanism was programmed in the solver by integrating three kinetic schemes: combustion, NOx formation, and NOx reduction. Regarding NOx reduction, the model of Miller and Glarborg [35], based on 24 species and 134 reactions, was employed. This model was validated with experimental results elsewhere [36], verifying that it reproduces the experimental trends with a reasonable accuracy. Regarding NOx formation, the so-called extended Zeldovich mechanism is usually employed in CFD to characterize NOx formation. This model is based on seven species and three reactions [37,38], listed in Appendix A. Since the goal of the present work was to study NOx , several NOx formation mechanisms were developed and validated with experimental results using 100% diesel fuel. Gasboard-3000 (Wuhan Cubic) series gas analyzers were employed to characterize the composition of the exhaust.

(6) experimental results elsewhere [36], verifying that it reproduces the experimental trends with a reasonable accuracy. Regarding NOx formation, the so-called extended Zeldovich mechanism is usually employed in CFD to characterize NOx formation. This model is based on seven species and three reactions [37,38], listed in Appendix A. Since the goal of the present work was to study NOx, several NOx formation mechanisms were Energies 2019, 12, 1255 6 of 13 developed and validated with experimental results using 100% diesel fuel. Gasboard-3000 (Wuhan Cubic) series gas analyzers were employed to characterize the composition of the exhaust gas, particularly Gasboard-3030 for HC andand Gasboard-3000 forfor NO, CO, gas, particularly Gasboard-3030 for HC Gasboard-3000 NO, CO,and andCO CO2.2 .This This gas gas analyzer provides numerical model and numerous numerous experimental experimental analyses analyses [39] provides NO instead of NOxx, but the numerical indicated x x is is primarily NO. TheThe extended Zeldovich model was compared to those Mellor indicated that thatNO NO primarily NO. extended Zeldovich model was compared toofthose of et al. [40], based six reactions and eight species; Kilpinen [41], based on based 10 reactions Mellor et al. [40],on based on six reactions and eight Zabetta species; and Zabetta and Kilpinen [41], on 10 and 11 species; Yangand et al. [42], based 20 species and 43and reactions, listed listed in Appendix B. The reactions and 11and species; Yang et al. [42],on based on 20 species 43 reactions, in Appendix B. results obtained by these kinetic models at several loadsloads and 1500 in Figure 4. This The results obtained by these kinetic models at several and rpm 1500 are rpmshown are shown in Figure 4. figure indicates that that these models provide a satisfactory This figure indicates these models provide a satisfactorycorrespondence correspondencebetween betweennumerical numerical and experimental NOxx trends. trends. Particularly, Particularly, the the extended extendedZeldovich Zeldovichmodel modelprovides providesaa13.9% 13.9%average averageerror, error, while Mellor [41] 8.1%, and Yang et et al.al. [42] 5.8%. TheThe R2 Mellor et etal. al.[40] [40]achieve achieve9.2%, 9.2%,Zabetta Zabettaand andKilpinen Kilpinen [41] 8.1%, and Yang [42] 5.8%. 2 resulted in 0.955, 0.958, 0.959, and 0.962 forforthe R resulted in 0.955, 0.958, 0.959, and 0.962 theextended extendedZeldovich, Zeldovich,Mellor Melloretetal. al. [40], [40], Zabetta Zabetta and Kilpinen [41], and Yang et al. [42] models, respectively. According to this, the model of Yang et al. [42] and Yang was used for for the the computations computationsin inthe theremainder remainderofofthe thepresent presentpaper. paper.. Figure 4. 4. NO NOxx emissions emissions experimentally experimentally and and numerically numerically obtained. obtained. Figure. Regarding the combustion mechanism, the simplest procedure is to assume that the kinetics are so Regarding the combustion mechanism, the simplest procedure is to assume that the kinetics are fast that chemical species remain at equilibrium [30,36,43]. This hypothesis is commonly used in CFD so fast that chemical species remain at equilibrium [30,36,43]. This hypothesis is commonly used in taking into account the high temperatures reached during combustion, which promote fast kinetics. CFD taking into account the high temperatures reached during combustion, which promote fast Nevertheless, a kinetic scheme is more accurate than the equilibrium hypothesis since the warming kinetics. Nevertheless, a kinetic scheme is more accurate than the equilibrium hypothesis since the in the expansion process of an engine and dilution with the excess air elongate the time needed to warming in the expansion process of an engine and dilution with the excess air elongate the time achieve equilibrium. For instance, in the case of CO and HC, one of the sources of formation are needed to achieve equilibrium. For instance, in the case of CO and HC, one of the sources of warm regions that are not able to burn properly. In this case, the chemical equilibrium hypothesis formation are warm regions that are not able to burn properly. In this case, the chemical equilibrium leads to overestimating the levels of these species. For this reason, a chemical kinetic model was hypothesis leads to overestimating the levels of these species. For this reason, a chemical kinetic employed in the present work. Particularly, Ra and Reitz’s [44] model, based on 131 reactions and model was employed in the present work. Particularly, Ra and Reitz’s [44] model, based on 131 41 species, was chosen. HC, CO, and CO emissions as well as brake-specific fuel consumption reactions and 41 species, was chosen. HC,2 CO, and CO2 emissions as well as brake-specific fuel at different loads and 1500 rpm are represented in Figure 5. This figure shows that a reasonable agreement was obtained, since the average error was 6.2%, 6.9%, 4.1%, and 3.8% for HC, CO, CO2 , and brake-specific fuel consumption (BSFC), respectively. Several reasons are responsible for the discrepancies between the numerical and experimental results obtained in Figures 4 and 5. Numerical techniques introduce inevitable errors due to the hypotheses assumed and the discretization on both the domain and governing equations. The RANS equations are time-averaged equations and the k-ε is a turbulence model that reproduces, as accurately as possible, the fluid flow. The P1 model for radiation, Kelvin-Helmholtz and Rayleigh-Taylor models for breakup, and Dukowicz model for heat-up and evaporation of the droplets are also models which reproduce, as accurately as possible, the.

(7) discretization on both the domain and governing equations. The RANS equations are time-averaged equations and the k-ε is a turbulence model that reproduces, as accurately as possible, the fluid flow. The P1 model for radiation, Kelvin-Helmholtz and Rayleigh-Taylor models for breakup, and Dukowicz model for heat-up and evaporation of the droplets are also models which reproduce, as accurately as possible, the physical mechanisms. Besides, the kinetic models employed are Energies 2019, 12, 1255 7 of 13 simplifications of the tens of species and hundreds of reactions involved in the chemistry of combustion. More accurate turbulence models such as LES (Large Eddy Simulation) or DNS (Direct Numerical Simulation)Besides, are too the computationally the simulationofofthe thetens hundreds of physical mechanisms. kinetic modelsexpensive, employedas areissimplifications of species reactions involved. With the advancement of computational resources, it is expected that in the near and hundreds of reactions involved in the chemistry of combustion. More accurate turbulence models future will(Large be possible to employ more accurate and more chemical reactions such asitLES Eddy Simulation) or DNS (Directturbulence Numericalmodels Simulation) are too computationally within a reasonable expensive, as is thecomputational simulation of time. the hundreds of reactions involved. With the advancement of Regarding resources, the simulation of fuel droplet Kelvin-Helmholtz and to Rayleigh-Taylor computational it is expected that inbreakup, the nearthe future it will be possible employ more breakup [33]models were employed, the heat-up andwithin evaporation of the droplets were modeled accurate models turbulence and moreand chemical reactions a reasonable computational time. by the Dukowicz model [34].. Figure 5.5.HC, CO,CO, and CO as well as fuel consumption experimentally Figure HC, and2 emissions CO2 emissions asbrake-specific well as brake-specific fuel (BSFC) consumption (BSFC) and numericallyand obtained. experimentally numerically obtained.. Regarding the simulation of fuel droplet breakup, the Kelvin-Helmholtz and Rayleigh-Taylor 3. Results and Discussion breakup models [33] were employed, and the heat-up and evaporation of the droplets were modeled Once the numerical model was validated, it offered an efficient, cheap, and fast method to by the Dukowicz model [34]. analyze the performance and emissions of the engine. In the present work, the model was employed to analyze and several injection rate shapes, particularly rectangular, triangular, and parabolic. Before 3. Results Discussion that, it was necessary to determine a proper NH3/NOi volumetric ratio, where NOi is the NO Once the numerical model was validated, it offered an efficient, cheap, and fast method to analyze concentration in the exhaust gas without ammonia injection. Figure 6 represents the NOx reduction the performance and emissions of the engine. In the present work, the model was employed to analyze as well as the ammonia slip in the exhaust gas against the NH3/NOi ratio using a rectangular several injection rate shapes, particularly rectangular, triangular, and parabolic. Before that, it was ammonia injection shape, start of ammonia injection 43.2º after top dead center at the end of necessary to determine a proper NH3 /NOi volumetric ratio, where NOi is the NO concentration in compression stroke, and 10º injection duration. This figure shows that more NOx reduction is the exhaust gas without ammonia injection. Figure 6 represents the NOx reduction as well as the obtained as NH3 is increased. Due to the toxicity of ammonia, it is important to maintain low values ammonia slip in the exhaust gas against the NH3 /NOi ratio using a rectangular ammonia injection of un-reacted ammonia slip to the exhaust. For this reason, a value of 2 was employed as a threshold shape, start of ammonia injection 43.2º after top dead center at the end of compression stroke, and 10º for NH3/NOi in the remaining discussion of the present paper. Values between 2 and 3 are injection duration. This figure shows that more NOx reduction is obtained as NH3 is increased. Due considered appropriate as well, since NOx reduction levels off around 3. A value greater than 3 to the toxicity of ammonia, it is important to maintain low values of un-reacted ammonia slip to the provides few improvements in NOx reduction but a considerably higher amount of un-reacted exhaust. For this reason, a value of 2 was employed as a threshold for NH3 /NOi in the remaining discussion of the present paper. Values between 2 and 3 are considered appropriate as well, since NOx reduction levels off around 3. A value greater than 3 provides few improvements in NOx reduction but a considerably higher amount of un-reacted ammonia slips into the atmosphere, since ammonia slip analyses showed a noticeable increment with an increasing NH3 /NOi ratio..

(8) Energies 2019, 2019, 12, 12, xx FOR FOR PEER PEER REVIEW REVIEW Energies. of 14 14 88 of. ammonia slips into Energies 2019,slips 12, 1255 ammonia into. the atmosphere, atmosphere, since since ammonia ammonia slip slip analyses analyses showed showed aa noticeable noticeable increment increment 8 of 13 the with an increasing NH 3/NOi ratio. with an increasing NH3/NOi ratio.. Figure 6. NOx reduction against the NH3 /NOi ratio. Start of ammonia injection 43.2º, 10º injection Figure 6. 6. NO NOxx reduction reduction against against the the NH NH33/NO /NOii ratio. ratio. Start Start of of ammonia ammonia injection injection 43.2º 43.2º,, 10º 10º injection injection Figure duration, rectangular injection shape. duration, rectangular rectangular injection injection shape. shape. duration,. Once the quantity of ammonia was determined, another important parameter to study was the Once the the quantity quantity of of ammonia ammonia was was determined, determined, another another important parameter parameter to to study study was was the the shapeOnce of ammonia injection. The injection shapes indicated important in Figure 7 were analyzed. These are shape of ammonia injection. The injection shapes indicated in Figure 7 were analyzed. These are shape of ammonia injection. The injection indicated in Figure 7 were analyzed. These are rectangular, triangular, and parabolic. Theseshapes three injection profiles refered to NH 3 /NOi = 2 and 10º rectangular, triangular, triangular, and and parabolic. parabolic. These These three three injection injection profiles profiles refered to to NH3/NOi = 2 and 10º rectangular, injection duration. At 1500 rpm, 10º injection duration corresponded refered to 0.0011 s.NH3/NOi = 2 and 10º injection duration. At 1500 rpm, 10º injection duration corresponded to 0.0011 s. injection duration. At 1500 rpm, 10º injection duration corresponded to 0.0011 s.. Figure 7. Ammonia injection rates analyzed. Figure 7. 7. Ammonia Ammonia injection injection rates rates analyzed. analyzed. Figure. The results for these three shapes are indicated in Figure 8, which illustrates the NOx reduction Thethe results for these three threeinjection shapes are are indicated in Figure Figure 8, which whichThis illustrates the NOxx reduction reduction against start for of ammonia using a 10º injection duration. figurethe indicates that the The results these shapes indicated in 8, illustrates NO against the start of ammonia injection using a 10º injection duration. This figure indicates that the the instant of ammonia injection has a using noticeable on the NOx reduction. Theindicates reason is that against start the start of ammonia injection a 10ºeffect injection duration. This figure instant start start ofx ammonia ammonia injection has aa noticeable noticeable effectin ona the the NOxxtemperature reduction. The The reason is that that the the process of NO reduction using ammonia is only efficient narrow range. Consequently, instant of injection has effect on NO reduction. reason is process of NO x reduction using ammonia is only efficient in a narrow temperature range. in an internal combustion engine, is too critical determine of ammonia injection due to process of NO x reduction usingit ammonia is to only efficienttheinstart a narrow temperature range. Consequently, in anin-cylinder internal combustion combustion engine, it is is too critical critical to determine determine the start start of of ammonia the variation of in thean temperatureengine, with the crankshaft angle. For the parameters analyzed in Consequently, internal it too to the ammonia injection due to the variation of the in-cylinder temperature with the crankshaft angle. For the Figure 8, the maximum NO reduction was obtained at start of ammonia injections 43.2º, 45.8º, and x injection due to the variation of the in-cylinder temperature with the crankshaft angle. For the parameters analyzed in Figure Figure 8, theofmaximum maximum NOstroke x reduction was obtained obtained at start startand of ammonia ammonia 42.9º after top dead center at the8, end compression for rectangular, triangular, parabolic parameters analyzed in the NO x reduction was at of injectionsshapes, 43.2º,, 45.8º 45.8º,, and and 42.9º 42.9º after top dead dead NO center at the the end end of68.3%, compression stroke injection respectively. The after corresponding wereof 66.8, andstroke 71.1% for x reductions injections 43.2º top center at compression rectangular, triangular, triangular,and and parabolic injection shapes, respectively. The which corresponding NOxx rectangular, parabolic injection shapes, respectively. The reason explains these rectangular, triangular, and parabolic injection shapes, respectively. The corresponding NO reductions were 68.3%, 66.8, and 71.1% for rectangular, triangular, and parabolic injection shapes, differences is that68.3%, high temperatures promote ammonia triangular, to oxidize to NO instead injection of reacting with reductions were 66.8, and 71.1% for rectangular, and parabolic shapes, respectively. The reason which explains these differences is that high temperatures promote NO, while at lower temperatures, the reduction reactions become slow and un-reacted ammonia can respectively. The reason which explains these differences is that high temperatures promote ammonia to oxidize to NO instead of reacting with NO, while at lower temperatures, the reduction be emitted. to this, the ammonia injected progressively from 0 tothe a maximum ammonia to According oxidize to NO instead of reactingmust with be NO, while at lower temperatures, reduction reactions become slow and and un-reacted ammoniashape can be be emitted. According to this,properly the ammonia ammonia value and become then decrease to 0 again. The parabolic is the shape According that adapts to more to this reactions slow un-reacted ammonia can emitted. this, the must be injected progressively from 0 to a maximum value and then decrease to 0 again. The tendency while the triangular shape is the most different. Regarding brake-specific consumption and must be injected progressively from 0 to a maximum value and then decrease to 0 again. The parabolic shape is the shape that adapts more properly to this tendency while the triangular shape is other emissions, such CO,that HC,adapts and CO verified thattendency the effectwhile was negligible. parabolic shape is the as shape more properly to this the triangular shape is 2 , it was.

(9) Energies 2019, 12, x FOR PEER REVIEW. 9 of 14. Energies 2019, 12, 1255. 13 the most different. Regarding brake-specific consumption and other emissions, such as CO, HC,9 of and CO2, it was verified that the effect was negligible.. Figure 8. NOx reduction against the start of ammonia injection for rectangular (green color), triangular Figure 8. NO reduction(blue against start of ammonia forduration. rectangular (green color), (red color), andx parabolic color)the injection shapes using injection 10º injection triangular (red color), and parabolic (blue color) injection shapes using 10º injection duration.. Another interesting parameter to analyze is the injection duration. Figure 9 shows the NOx Another interesting parameter to analyze is the injection Figuredata, 9 shows the NOx reduction against the duration of ammonia injection. In order duration. to obtain these the ammonia reduction against the duration of ammonia injection. In order to obtain these data, the ammonia injection injection was equally shortened at start and end, using the same quantity of ammonia injected. The was equallyNO shortened at start and end, the same quantitytoofthe ammonia injected. maximum obtained wasusing 78.2%, corresponding parabolic shapeThe at 1ºmaximum injection x reduction NOx reduction obtained was to the at 1ºwas injection duration. can duration. As can be seen in 78.2%, Figure corresponding 9, as the duration ofparabolic ammoniashape injection shortened, theAsNO x be seen in Figure 9, as the duration of ammonia injection was shortened, the NO x reduction increased. reduction increased. The reason is that lower injection durations facilitate the supply of ammonia The reason is thatcrankshaft lower injection the supplyin-cylinder of ammonia at the optimum at the optimum angledurations and thusfacilitate at the optimum temperature. On crankshaft the other angle longer and thus at the optimum in-cylinder temperature. On the other hand, injections deviate hand, injections deviate more from the optimum crankshaft angle and longer thus from the optimum more from the optimum crankshaft angle and thus fromfrom the optimum in-cylinder temperature in-cylinder temperature to reduce NO the data obtained from Figure 9, to thereduce NOx x . Particularly, NOx. Particularly, from the 66.8% data obtained from Figure 9, thewith NOx10º reduction improved from 66.8%for forthe the reduction improved from for the triangular shape injection duration to 78.2% triangular shape with 10º injection duration to 78.2% for the parabolic shape with 1º injection duration; parabolic shape 1º injection duration; i.e., a 11.4% NOx improvement was obtained. Another i.e., a 11.4%that NOxcan improvement obtained. conclusion thatbetween can be obtained from triangular, Figure 9 is conclusion be obtainedwas from Figure 9Another is that the differences rectangular, that parabolic the differences rectangular, and parabolic shapes becomeis smaller as the and shapes between become smaller as the triangular, injection duration is reduced. The reason that at shorter injection duration is reduced. The reason is that at shorter injection rates the ammonia is only injected at injection rates the ammonia is only injected at the optimum crankshaft angle and the injection duration the and the injection duration is so applications, short that thethe shape has arate negligible is sooptimum short thatcrankshaft the shape angle has a negligible influence. In practical injection can be influence. In practical applications, the injection rate can be shortened to the technically feasible level. shortened to the technically feasible level. Regarding brake-specific consumption and other emissions, Energies 2019, 12, x FOR PEER REVIEW 10 of 14 Regarding brake-specific emissions, such CO, HC, and CO2, it was verified that such CO, HC, and CO2 , itconsumption was verifiedand thatother the effect was negligible. the effect was negligible.. Figure 9. NOx reduction against the duration of injection for rectangular (green color), triangular (red color),Figure and parabolic (blue color) injection shapes. 9. NOx reduction against the duration of injection for rectangular (green color), triangular (red color), and parabolic (blue color) injection shapes.. 4. Conclusions This paper analyzes the capability of ammonia to reduce NOx in a hydrogen-diesel engine. The motivation comes from the importance of the reduction of NOx emissions which are usually produced.

(10) Energies 2019, 12, 1255. 10 of 13. 4. Conclusions This paper analyzes the capability of ammonia to reduce NOx in a hydrogen-diesel engine. The motivation comes from the importance of the reduction of NOx emissions which are usually produced when hydrogen is added to internal combustion engines, and the increasingly stringent NOx restrictions. A CFD model was developed to analyze the commercial engine MAN D2840LE V10. Since ammonia is toxic, it is important to optimize NOx reduction, minimizing the un-reacted ammonia slipping into the atmosphere. According to this, a NH3 /NOi proportion around 2 was considered adequate. It was verified that the NOx reduction process using ammonia is highly dependent on the in-cylinder temperature. According to this, several ammonia injection shapes (rectangular, triangular, and parabolic) and durations (from 1º to 10º) were analyzed. The most appropriate case provided a 78.2% NOx reduction using a parabolic injection shape and 1º injection duration. On the other hand, the worst case analyzed provided a 66.8% NOx reduction using a triangular shape and 10º injection duration. As the injection duration is reduced, the differences between rectangular, triangular, and parabolic shapes become practically negligible. In practical applications, the injection rate can be shortened to the level that is technically feasible. For instance, in the present engine, which runs at 1500 rpm, 1º crankshaft angle corresponds to 0.00011 s. Current electromagnetic or solenoid injectors have longer response times but some piezo-injectors are able to switch on an off in tens of microseconds. These short injections provide practically identical results for rectangular, triangular, and parabolic injection shapes. The main contribution of the present paper is that it provides a cheap and fast tool for studying NOx reduction in internal combustion engines. This tool constitutes an alternative to expensive and laborious experimental tests. It can be employed to study the influence of parameters such as the exhaust and intake pressures, number and position of injectors, nozzles, injection pressure, cam profile design, compression ratio, etc. In future works, the goal is to employ this numerical model to study other engine operation conditions and other emission reduction procedures, such as water addition, common rail, exhaust gas recirculation, etc. Furthermore, further important work is needed to develop more tests in this and other engines in order to obtain a more complete validation of the numerical model. Author Contributions: Conceptualization, M.I. Lamas and C.G. Rodriguez; methodology, M.I. Lamas and C.G. Rodriguez; software, M.I. Lamas and C.G. Rodriguez; validation, M.I. Lamas and C.G. Rodriguez; formal analysis, M.I. Lamas and C.G. Rodriguez; investigation, M.I. Lamas and C.G. Rodriguez; resources, M.I. Lamas and C.G. Rodriguez; writing—original draft preparation, M.I. Lamas; writing—review and editing, M.I. Lamas. Funding: This research received no external funding. Acknowledgments: The authors would like to express their gratitude to Talleres Pineiro S.L., sale and repair of marine engines, as well as to Norplan Engineering S.L., and recommend the courses on CFD with OpenFOAM and C++ applied to OpenFOAM available at www.technicalcourses.net. Conflicts of Interest: The authors declare no conflict of interest.. Appendix A Reaction equations of the Zeldovich mechanism. The rate coefficients are in the form k f = AT b e−E0 /T (Table A1). Table A1. Zeldovich mechanism.. 1 2 3. Reaction. A (cm3 /mol s). b. E0 (K). N2 + O ↔ NO + N N + O2 ↔ NO + O N + OH ↔ NO + H. 1.800·1014 1.800·1010 7.100·1013. 0 1 0. 38370 4680 450.

(11) Energies 2019, 12, 1255. 11 of 13. Appendix B Reaction equations of the Yang et al. [42] mechanism. The rate coefficients are in the form k f = AT b e−E0 /T (Table A2). Table A2. Yang et al. [42] mechanism.. 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43. Reaction. A (cm3 /mol s). b. E0 (K). N2 + O ↔ NO + N N + O2 ↔ NO + O N + OH ↔ NO + H N2 O + O ↔ 2NO N2 O + O ↔ N2 + O2 N2 O + H ↔ N2 + OH N2 O + N2 ↔ 2N2 + O N2 O + N ↔ NO + N2 N2 O + NO ↔ NO2 + N2 N2 O + O2 ↔ NO + NO2 NO2 + O ↔ NO + O2 NO2 + H ↔ NO + OH NO2 + NO2 ↔ 2NO + O2 NO2 + N ↔ O + N2 O NO + HO2 ↔ NO2 + OH NO + N2 + O2 ↔ N2 O + NO2 N2 + O2 ↔ 2NO NO + O2 ↔ N + O3 NO2 + O ↔ N + O3 NO + O3 ↔ O + NO3 N2 + HO2 ↔ NO + HNO N + OH ↔ NH + O N + H2 ↔ NH + H N + H2 O ↔ NH + OH NH + OH ↔ NO + H2 NH + O ↔ NO + H NH + OH ↔ HNO + H NH + O2 ↔ HNO + O HNO ↔ H + NO HNO + OH ↔ NO + H2 O HNO + H ↔ NO + H2 HNO + O ↔ NO + OH NH + NO ↔ N2 O + H N2 O + NH ↔ N2 + HNO NO2 + NH ↔ HNO + NO N + H2 O ↔ NH + OH N + HO2 ↔ NH + O2 NO + HO2 ↔ HNO + O HNO + NO ↔ N2 O + OH HNO + N ↔ NO + NH HNO + HNO ↔ 2NO + H2 NH + N ↔ N2 + H NH + N2 ↔ N + H + N2. 1.800·1014. 0.0 1.0 0.0 0.0 0.0 0.0 –2.5 0.0 0.0 –1.5 0.0 0.00 0.00 0.0 0.0 0.0 0.0 –1.0 –0.5 –1.5 0.5 0.0 0.0 0.0 0.56 0.5 0.0 0.0 0.0 0.5 0.0 0.5 0.0 0.0 0.5 0.0 0.0 0.0 0.0 0.0 0.5 0.5 –2.0. 38370 4680 450 13400 14200 7600 32170 10000 25000 4985 300 740 13500 0 –240 18000 53800 63140 40280 8000 21550 2165 11230 18430 755 0.0 1460 6546 25179 0 1510 0 10600 3000 2000 1000 1000 1000 17100 1000 2230 0 42000. 1.800·1010 7.100·1013 6.900·1013 1.000·1014 7.587·1013 6.918·1023 1.000·1013 1.000·1014 6.000·1014 1.000·1013 3.467·1014 1.995·1012 5.012·1012 2.089·1012 2.300·1014 2.730·1013 2.700·1014 3.700·1014 6.000·1014 5.925·1010 1.290·1014 1.320·1015 3.590·1015 1.600·1012 5.000·1011 6.440·1011 4.380·1012 1.900·1016 2.100·1012 1.400·1013 5.000·1011 2.240·1013 1.995·1012 1.000·1011 1.000·1013 1.000·1013 1.900·1011 6.140·1012 1.000·1013 2.000·1010 6.310·1011 3.160·1021. References 1. 2.. Seddiek, I.S.; Elgohary, M.M.; Ammar, N.R. The hydrogen-fuelled internal combustion engines for marine applications with a case study. Brodogradnja 2015, 66, 23–38. Lamas, M.I.; Rodríguez, C.G.; Rodríguez, J.D.; Telmo, J. Internal modifications to reduce pollutant emissions from marine engines. A numerical approach. J. Nav. Archit. Mar. Eng. 2013, 5, 493–501. [CrossRef].

(12) Energies 2019, 12, 1255. 3. 4.. 5. 6. 7.. 8.. 9. 10.. 11.. 12. 13. 14. 15. 16. 17. 18. 19.. 20. 21. 22. 23.. 24.. 12 of 13. Seddiek, I.S.; Elgohary, M.M. Eco-friendly selection of ship emissions reduction strategies with emphasis on SOx and NOx emissions. Int. J. Nav. Arch. Ocean Eng. 2014, 6, 737–748. [CrossRef] Ovaska, T.; Niemi, S.; Sirvio, K.; Nilson, O.; Portin, K.; Asplund, T. Effects of alternative marine diesel fuels on the exhaust particle size distributions of an off-road diesel engine. Appl. Therm. Eng. 2019, 150, 1168–1176. [CrossRef] Abdulwahid, A.A.; Sigu, R.; Brown, R.J. Underground diesel exhaust wet scrubbers: Current status and future prospects. Energies 2018, 11, 3006. [CrossRef] Frigo, S.; Gentili, R. Analysis of the behavior of a 4-stroke SI engine fuelled with ammonia and hydrogen. Int. J. Hydrogen Energy 2013, 38, 1607–1615. [CrossRef] Reyes-Ramírez, I.; Martínez-Boggio, S.D.; Curto-Risso, P.L.; Medina, A.; Calvo Hernández, A.; Guzmán-Vargas, L. Symbolic analysis of the cycle-to-cycle variability of a gasoline–hydrogen fueled spark engine model. Energies 2018, 11, 968. [CrossRef] Barrios, C.C.; Domínguez-Sáez, A.; Hormigo, D. Influence of hydrogen addition on combustion characteristics and particle number and size distribution emissions of a TDI diesel engine. Fuel 2017, 199, 162–168. [CrossRef] Li, Y.; Chen, H.; Zhang, X.; Tan, C.; Ding, Y. Renewable energy carriers: Hydrogen or liquid air/nitrogen? Appl. Therm. Eng. 2010, 30, 1985–1990. [CrossRef] Wang, W.; Herreros, J.M.; Tsolakis, A.; York, A.P.E. Ammonia as hydrogen carrier for transportation; investigation of the ammonia exhaust gas fuel reforming. Int. J. Hydrogen Energy 2013, 38, 9907–9917. [CrossRef] Cervantes-Bobadilla, M.; Escobar-Jiménez, R.F.; Gómez-Aguilar, J.F.; García-Morales, J.; Olivares-Peregrino, V.H. Experimental study on the performance of controllers for the hydrogen gas production demanded by an internal combustion engine. Energies 2018, 11, 2157. [CrossRef] Boretti, A. Novel dual fuel diesel-ammonia combustion system in advanced TDI engines. Int. J. Hydrogen Energy 2017, 42, 7071–7076. [CrossRef] Comotti, M.; Frigo, S. Hydrogen generation system for ammonia-hydrogen fuelled internal combustion engines. Int. J. Hydrogen Energy 2015, 40, 10673–10686. [CrossRef] Lee, J.H.; Kim, J.H.; Park, J.H.; Kwon, O.C. Studies on properties of laminar premixed hydrogen-added ammonia/air flames for hydrogen production. Int. J. Hydrogen Energy 2010, 35, 1054–1064. [CrossRef] Ryu, K.; Zacharakis-Jutz, G.E.; Kong, S.C. Performance enhancement of ammonia-fueled engine by using dissociation catalyst for hydrogen generation. Int. J. Hydrogen Energy 2014, 39, 2390–2398. [CrossRef] Boretti, A.A. Novel heavy duty engine concept for operation dual fuel H2-NH3. Int. J. Hydrogen Energy 2012, 37, 7869–7876. [CrossRef] Sahin, Z.; Akcanca, I.Z.; Durgun, O. Experimental investigation of the effects of ammonia solution (NH3OH) on engine performance and exhaust emissions of a small diesel engine. Fuel 2018, 214, 330–341. [CrossRef] Reiter, A.J.; Kong, S.C. Combustion and emissions characteristics of compression-ignition engine using dual ammonia-diesel fuel. Fuel 2011, 90, 87–97. [CrossRef] Gill, S.S.; Chatha, G.S.; Tsolakis, A.; Golunski, S.E.; York, A.P.E. Assessing the effects of partially decarbonizing a diesel engine by co-fuelling with dissociated ammonia. Int. J. Hydrogen Energy 2012, 37, 6074–6083. [CrossRef] Nishi, M.; Kanehara, M.; Lida, N. Assessment for innovative combustion on HCCI engine by controlling EGR ratio and engine speed. Appl. Therm. Eng. 2016, 99, 42–60. [CrossRef] Wu, H.W.; Hsu, T.T.; He, J.Y.; Fan, C.M. optimal performance and emissions of diesel/hydrogen-rich gas engine varying intake air temperature and EGR ratio. Appl. Therm. Eng. 2017, 124, 381–392. [CrossRef] di Sarli, V. Stability and emissions of a lean pre-mixed combustor with rich catalytic/lean-burn pilot. Int. J. Chem. React. Eng. 2014, 12, 77–89. [CrossRef] Lamas, M.I.; Rodríguez, C.G.; Aas, H.P. Computational fluid dynamics analysis of NOx and other pollutants in the MAN B&W 7S50MC marine engine and effect of EGR and water addition. Int. J. Marit. Eng. 2013, 155, A81–A88. Finesso, R.; Hardy, T.; Mancarella, A.; Marello, O.; Mittica, A.; Spessa, E. Real-time simulation of torque and nitrogen oxide emissions in a 11.0 L heavy-duty diesel engine for model-based combustion control. Energies 2019, 12, 460. [CrossRef].

(13) Energies 2019, 12, 1255. 25. 26.. 27. 28. 29. 30.. 31. 32. 33. 34. 35. 36. 37. 38. 39. 40. 41.. 42.. 43. 44.. 13 of 13. Jasminska, N.; Brestovic, T.; Puskar, M.; Grega, R.; Rajzinger, J.; Korba, J. Evaluation of hydrogen storage capacities on individual adsorbents. Meas. J. Int. Meas. Confed. 2014, 56, 219–230. [CrossRef] Lamas, M.I.; Rodríguez, J.d.; Castro-Santos, L.; Carral, L.M. Effect of multiple injection strategies on emissions and performance in the Wartsila 6L 46 marine engine. A numerical approach. J. Clean. Prod. 2019, 206, 1–10. [CrossRef] Park, C.; Park, W.; Kim, Y.; Choi, Y.; Lim, B. Effect of valve timing and excess air ratio on torque in hydrogen-fueled internal combustion engine for UAV. Energies 2019, 12, 771. [CrossRef] Miyamoto, N.; Ogawa, H.; Wang, J.; Shudo, T. Yamazaki K. Diesel NOx reduction with ammonium deoxidizing agents directly injected into the cylinder. Int. J. Veh. Des. 1995, 16, 71–79. [CrossRef] Larbi, N.; Bessrour, J. Measurement and simulation of pollutant emissions from marine diesel combustion engine and their reduction by ammonia injection. Adv. Mech. Eng. 2010, 41, 898–906. [CrossRef] Lamas, M.I.; Rodriguez, C.G.; Rodriguez, J.D.; Telmo, J. Computational fluid dynamics of NOx reduction by ammonia injection in the MAN B&W 7S50MC marine engine. Int. J. Marit. Eng. 2014, 156, A213–A220. [CrossRef] Bosse, T.K. High Temperature Gas Dynamics, 1st ed.; Springer: Berlin/Heidelberg, Germany; New York, NY, USA, 2004. Lamas, M.I.; Rodríguez, C.G. CFD analysis of the scavenging process in the MAN B&W 7S50MC two-stroke diesel marine engine. J. Ship Res. 2012, 56, 154–161. [CrossRef] Ricart, L.M.; Xin, J.; Bower, G.R.; Reitz, R.D. In-Cylinder Measurement and Modeling of Liquid Fuel Spray Penetration in a Heavy-Duty Diesel Engine; SAE 971591; SAE International: Warrendale, PA, USA, 1997. Dukowicz, J.K. Quasi-Steady Droplet Change in the Presence of Convection; LA7997-MS; Informal Report Los Alamos Scientific Laboratory: Los Alamos, NM, USA, 1979. Miller, J.A.; Glarborg, P. Modeling the formation of N2O and NO2 in the thermal DeNOx process. Springer Ser. Chem. Phys. 1996, 61, 318–333. Lamas, M.; Rodriguez, C.G. Numerical model to analyze NOx reduction by ammonia injection in diesel-hydrogen engines. Int. J. Hydrogen Energy 2017, 42, 26132–26141. [CrossRef] Zeldovich, Y.B.; Sadovnikov, D.A.; Kamenetskii, F. Oxidation of Nitrogen in Combustion; Institute of Chemical Physics: Moscow-Leningrad, Russia, 1947. Lavoie, G.A.; Heywood, J.B.; Keck, J.C. Experimental and theoretical investigation of nitric oxide formation in internal combustion engines. Combust. Sci. Technol. 1970, 1, 313–326. [CrossRef] Woodyard, D. Pounder’s Marine Diesel Engines and Gas Turbines, 8th ed.; Elsevier: Oxford, UK, 2004; ISBN 9780080943619. Mellor, A.M. Skeletal Mechanism for NOx Chemistry in Diesel Engines; SAE Tech Pap 981450; SAE International: Warrendale, PA, USA, 1998. Zabetta, E.C.; Kilpinen, P.; Hupa, M.; Stähl, K.; Leppälahti, J.; Cannon, M.; Nieminen, J. Kinetic modeling study on the potential of staged combustion in gas turbines for the reduction of nitrogen oxide emissions from biomass IGCC plants. Energy Fuels 2000, 14, 751–761. [CrossRef] Yang, H.; Krishnan, S.R.; Srinivasan, K.K.; Midkiff, K.C. Modeling of NOx emissions using a super-extended Zeldovich mechanism. In Proceedings of the ICEF03 2003 Fall Technical Conference of the ASME Internal Combustion Engine Division, Erie, PA, USA, 7–10 September 2003. Lamas, M.I.; Rodríguez, C.G. Numerical model to study the combustion process and emissions in the Wärtsilä 6L 46 four-stroke marine engine. Pol. Marit. Res. 2013, 20, 61–66. [CrossRef] Ra, Y.; Reitz, R. A reduced chemical kinetic model for IC engine combustion simulations with primary reference fuels. Combust. Flame 2008, 155, 713–738. [CrossRef] © 2019 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY) license (http://creativecommons.org/licenses/by/4.0/)..

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